Ultra-resilient multi-layer fluorinated diamond like carbon hydrophobic surfaces

Seventy percent of global electricity is generated by steam-cycle power plants. A hydrophobic condenser surface within these plants could boost overall cycle efficiency by 2%. In 2022, this enhancement equates to an additional electrical power generation of 1000 TWh annually, or 83% of the global solar electricity production. Furthermore, this efficiency increase reduces CO2 emissions by 460 million tons /year with a decreased use of 2 trillion gallons of cooling water per year. However, the main challenge with hydrophobic surfaces is their poor durability. Here, we show that solid microscale-thick fluorinated diamond-like carbon (F-DLC) possesses mechanical and thermal properties that ensure durability in moist, abrasive, and thermally harsh conditions. The F-DLC coating achieves this without relying on atmospheric interactions, infused lubricants, self-healing strategies, or sacrificial surface designs. Through tailored substrate adhesion and multilayer deposition, we develop a pinhole-free F-DLC coating with low surface energy and comparable Young’s modulus to metals. In a three-year steam condensation experiment, the F-DLC coating maintains hydrophobicity, resulting in sustained and improved dropwise condensation on multiple metallic substrates. Our findings provide a promising solution to hydrophobic material fragility and can enhance the sustainability of renewable and non-renewable energy sources.

showing that the 3-year durability tested FDLC (HTC3-yr tested) sample shows <30% reduction in condensation heat transfer coefficient compared to a freshly coated FDLC sample (HTCfresh FDLC). This is due to the reduction of surface wettability and increment of the condensate size imposing higher thermal resistance to heat transfer. However, the HTC3-yr tested is still ~74% higher than uncoated copper sample (HTCuncoated) which underwent filmwise condensation. Error bars were computed using the propagation of error (see Supplementary Note 4c). See Supplementary Note 7 for the details of the experiments.  S-12  Time-lapse photographs of durability tests during steam condensation on HTMS coated aluminum samples exposed to pure steam having (a) ~2 kPa vapor pressure and ~35℃ vapor temperature, and (b) ~25 kPa vapor pressure and ~65℃ vapor temperature. Surfaces exposed to high vapor pressure experience accelerated condensation rate, which results in much faster degradation of the hydrophobic chemistry as observed in (b). See Supplementary Note 9 for the details of the experiments. S-17 Supplementary Table 1. Summary of past studies on F-DLC conducted over the past 30 years. Studies are sorted focusing on studied substrates, fabricated layer thickness, number of layers, reported thermo-mechanical robustness and condensation durability. Here, Si = Silicon, SiC = Silicon Carbide, SiN= Silicon Nitride, SS = Stainless Steel, Cu = Copper, PMMA = Polymethyl Methacrylate, Al= Aluminum, KBr = Potassium bromide, PC = Polycabonate, PES = Polyethersulfone, Ti= Titanium. HTC= Heat Transfer Efficient, COF = Coefficient of Friction. Empty cells represent that the study did not report the data or did not conduct the experiments.

No
Year  Figure 1). The coating should also have low thermal resistance (<5 µm thick with high thermal conductivity) to condensation heat transfer and should sustain elevated thermo-mechanical environments and humid environments for long times (years) (Supplementary Figure 1).

Supplementary Note 2. Potential of DLC and Summary of Previous Studies
The first report on hard amorphous carbon films (also known as diamond like carbon, DLC) was published in 1953 by Heinz Schmellenmeier. About 20 years later, worldwide intensive research activities initiated on DLC. In the following years, the number of publications increased continuously and the importance for industrial applications became more and more evident.
Several deposition techniques were applied to prepare hydrogenated (a-C:H), non-hydrogenated Multi-layer DLC is more stable than monolayer DLC due to reduced stress. Inclusion of F also improves the flexibility of DLC, which is helpful in application where deformation, movement (tubes, guidewires) exists 37  covering scalable and substrate versatile fabrication methods using the design strategies for strong adhesion, thermo-mechanical robustness, and long-term durability to steam condensation with enhanced heat transfer performance.

Supplementary Note 3. Pathway from DLC to Multi-layer Durable F-DLC Development
Prior to developing the F-DLC formulation, we began by investigating the condensation  angles. However, the contact angle hysteresis is still high leading to undesirable condensation performance. To develop a low hysteresis and durable F-DLC coating we started to study the degradation mechanism and required properties to overcome the degradation, and finally, developed the multi-layer co-deposited F-DLC coating (Supplementary Figure 2ii).

Supplementary Note 4a. Chamber Setup
The complete description of the experimental setup (Supplementary Figure 3) with all components is discussed in detail in previous work. 124,125 In brief, the individual components are discussed below.
Main Chamber. The custom environmental chamber used for this work was designed by Gladwin Tank Manufacturing to withstand 5066 kPa of internal pressure and is ASME code stamped. The Vapor Generation. Water vapor is generated in a smaller stainless-steel vessel (8" OD Standard CF Tee, Kurt J. Lesker). The vessel is wrapped with three independently controlled 120 W tape heaters and is insulated to limit losses to the environment. The temperature of the fluid is monitored using a T-type thermocouple inserted directly into the liquid with a feedthrough connection on the top flange. For safety, a pressure relief valve set at 300 kPa is connected to the vapor generator.
Test Sample. The test samples used in this study are tubes with an outer diameter of 9.53 mm and a length of 134.6 mm. The diameter was chosen to maintain high heat transfer coefficients on the internal surface. A chilled water flow loop is used to cool the test samples to promote condensation on the F-DLC coated or bare copper surfaces. All auxiliary tubing inside the chamber is insulated to prevent condensate formation on its surface.

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Measurements and Data Acquisition (DAQ). The water flow rate was measured using an electromagnetic flow meter (Part #FMG93, Omega) and the inlet and outlet temperatures into the test section were measured using class A resistance temperature detectors, or RTDs (Part #AT-PX1123Y-LR4S1T2T, ReoTemp). During each test, the pressure of the working fluid (water) was measured using two pressure transducers (Baratron 728A and MicroPirani 925 supplied by MKS Instruments). A detailed reasoning for using data from the Baratron 728A over the classical MicroPirani 925 is listed in a previous work. 125 The vapor temperature and the chamber and test section wall temperatures are monitored with T-type thermocouples (Part #SCPSS-032, Omega).
All measurements were collected with a data acquisition system (DAQ) (PXIe 1073, National Instruments) and analyzed with LabVIEW. The PXIe system was specifically chosen due to its low DAQ error, enabling high fidelity measurements.
Finally, visual recordings of the condensation process were captured with a high-speed camera (Phantom v7.1, Vision Research) and a digital SLR camera (Pentax K-50).

Supplementary Note 4b. Experimental Procedure
The experimental procedure for steady-state vapor condensation within the chamber is discussed in detail in our previous work. 125 Here we discuss in brief the initial chamber preparation step and the steady state condensation operation.
A set of strict procedures were followed to maintain consistency across experiments and to ensure repeatable results. The test sample was connected to the coolant line using a T-joint Swagelok tube fitting and all connections were checked before the chamber was sealed to ensure there were no leaks. Once the environmental chamber front flange door was sealed, the vapor generator was completely filled with approximately 4.5 liters of deionized (DI) water and heated. After the DI water boiled (>95°C) for at least 10 minutes, the vacuum-pump-down procedure of the environmental chamber was initiated.
During the vacuum-pump-down, the pressure inside the chamber was monitored using both of the installed pressure transducers. This process took 30-45 minutes in order to achieve the target vacuum conditions. The environmental chamber was evacuated to a pressure of < 1 ± 0.025 Pa with a leak rate of 0.1 Pa.min -1 after chamber isolation. Evacuation eliminated noncondensable gases (NCGs) which add diffusional resistance to condensation heat transfer. 126 The final step in the startup procedure was to initiate the cooling water flow loop. The Neslab System 3 water pump reservoir was filled and turned on to the maximum recirculating pressure valve setting, giving a flow rate of ≈ 8 ± 0.2 L.min -1 . The relatively high flow rate of 8 L.min -1 was chosen to ensure a well-mixed highly turbulent flow with Reynolds number of Re d ID = w ̅ d ID / w = 24,000. This helped in maximizing the internal heat transfer coefficient and therefore reduce the overall condensation thermal resistance. The reservoir temperature was set at 7 ± 1°C and the coolant was circulated through the flow loop until the test section inlet temperature ( in) ranged between 7 to 8°C, and was maintained for at least 5 minutes to confirm steady state operation.
In summary, the following conditions were met before injecting vapor to the environmental chamber: 1) the liquid inside the vapor generator must reach the boiling point and generate vapor for at least 10 minutes, 2) the inlet coolant temperature must be between 7°C and 8°C for at least 5 minutes, and 3) the chamber pressure must be below 1 Pa. Once these conditions were reached, S-39 the valve in the vacuum-line was closed and the vacuum pump turned off. The vent valve on the vapor generator was closed and the valve connecting the vapor generator to the environmental chamber was gradually opened to allow vapor inside the chamber until it reached a desired pressure. All thermocouples, RTDs and pressure transducer readings were recorded over a 10minute period to obtain steady state condensation measurements.

Supplementary Note 4c. Heat Transfer Calculations and Error Analysis
Overall heat transfer coefficient ( ̅ ). The overall condensation heat transfer rate ( ) was calculated using an energy balance on the cooling water flowing inside the tube, as shown in Equation S1 : where is the overall condensation heat transfer rate, ̇ is the cooling water mass flow rate, p is the liquid cooling water specific heat, and out and in are the outlet and inlet temperatures, respectively. The overall heat transfer rate ( ) was then balanced with the overall heat transfer coefficient, ̅ as: where o is the tube outer surface area ( o = OD , where OD = 9.35 mm, = 134.62 mm) and ∆ LMTD is the log mean temperature difference defined by: 127 where v is the temperature of the surrounding saturated vapor inside the chamber ( v = sat ( v )).
The overall heat transfer coefficient ( ̅ ), which is only a function of experimentally obtained parameters, can thus be calculated as: Internal convection heat transfer coefficient (ℎ ). The calculated ̅ is a measure of the overall heat transfer performance from the vapor to the cooling water. It includes the convective resistances on the inner and outer walls and the conductive resistance through the copper wall and hydrophobic coating. Further calculations were performed to isolate the thermal resistance on the outer wall to quantify the condensation heat transfer coefficient, ℎ c , as measured from the vapor to the tube outer surface.
To extract ℎ c , the conductive resistance was calculated using the thermal conductivity of hollow conductive cylinder assembly and the internal resistance was calculated by estimating the internal heat transfer coefficient. The water-side heat transfer coefficient (ℎ i ) was estimated by the In Equations S5 to S7, represents the pipe friction factor, Re is the cooling water Reynolds number, Pr is the cooling water Prandtl number, is the cooling water density, i is the cooling water thermal conductivity, and b and s are the cooling water dynamic viscosities at the bulk and tube wall temperatures, respectively.

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Thermal resistances. The overall heat transfer performance from vapor-to-cooling water can be expressed as a series sum of four individual thermal resistances: the internal convective resistance conv , the radial conduction tube wall resistance wall , the conductive resistance through the coating, coating , and condensation heat transfer thermal resistance condensation , as shown in Considering the coating thickness of the F-DLC as h, the thermal resistance of the coating on a tube with length, L is: where OD is the outer diameter of the bare copper tube and F−DLC is the thermal conductivity of the F-DLC multilayer coating. Similarly, the conductive tube wall resistance can be calculated as: where OD and ID are the outer and inner bare tube diameters, respectively and Cu is the thermal conductivity of multipurpose copper. The internal convection resistance and external condensation resistance can be calculated as: (S11)
The overall heat transfer coefficient can be expressed as Equation S12 by considering all four thermal resistances in series: where o is the tube outer surface area ( o = OD ). Knowing ℎ i , a closed form solution can be obtained for ℎ c by combining all the relevant thermal resistances (internal convection and radial conduction through the tube wall and coating): (S13) Tube surface temperature ( surf ). The tube surface outer temperature, surf , was used to calculate the supersaturation for each test condition. The outer wall temperature was calculated using the total heat transfer rate and the conductive and water-side convective thermal resistances, as shown in Equation S14: where avg = ( out + in )/2. Rearranging Eq. S14, the tube surface temperature can be calculated as: Finally, the supersaturation, , defined as the ratio of the vapor pressure to the saturation pressure corresponding to the tube sample surface temperature is given by: . (S16)

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Error analysis. The uncertainty of the overall heat transfer coefficient, ̅ , was calculated by propagating the instrument uncertainty of each measured variable (Supplementary Table S3), as shown in Equation S17 : (S17) As the condensation heat transfer coefficient, ℎ c is a product of powers, the error is determined as a function of the first partial derivatives of ℎ with respect to its components: where t is the thermal resistance of the tube wall given by: S-45

Supplementary Note 4d. Dropwise Condensation Model
To calculate the overall condensation heat transfer, we first calculate the heat transfer through a water droplet. Then, we multiply the individual droplet heat transfer with the droplet distribution density to get the overall condensation heat flux. To calculate the droplet heat transfer, we studied the thermal resistance network. For a droplet residing on a surface, the dominating thermal resistances are the conduction resistance within the droplet, the surface conduction thermal resistance (coating), and the interfacial thermal resistance at the liquid-vapor interface. The conduction resistance is determined by following our previous study. 129 Due to the sufficiently small size of condensing droplets, convection inside of the droplet is low and is neglected. 130 Here we consider a droplet with radius R on a plain surface which is coated with F-DLC, as shown in Supplementary Figure 5. The contact angle of the F-DLC coated hydrophobic surface is assumed to be fixed regardless of the droplet size, surface temperature and vapor temperature. Knowing the vapor saturation temperature, sat and F-DLC coated tube surface temperature, surf , the heat transfer rate can be calculated by considering all thermal resistances.
Considering the droplet nucleation site density, Ns, 131 the effective coalescence radius of the droplets, Re above which droplets starts to coalesce can be calculated by: Droplet critical nucleation radius ( min ) was calculated by: 130 .
where , , and ℎ fg are the liquid-vapor surface tension, density, and latent heat of vaporization of water, respectively.

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The heat transfer rate through a single droplet with radius R is: where is the accommodation coefficient ranging from 0 to 1, representing the ratio of vapor molecules that will be captured by the liquid phase to the total number of vapor molecules reaching the liquid phase, and g is the specific gas constant of vapor. To The overall condensation heat transfer coefficient is calculated by: where Δ is the temperature difference between the ambient vapor saturation temperature, sat and the condensing surface temperature, surf (i.e. Δ = sat − surf ).

Supplementary Note 4e. Filmwise Condensation Heat Transfer Model
To model filmwise condensation of steam on the hydrophilic copper tube samples, the classical Nusselt model was used: 127,135 where is acceleration due to gravity ( = 9.81 m.s -2 ), v is the vapor density, l is the condensate liquid density, l is the condensate liquid dynamic viscosity, ℎ fg ′ is the modified latent heat of vaporization accounting for the change in specific heat of the condensate and p,l is the condensate liquid specific heat. S-49

Supplementary Note 5a. Test Facility
The durability tests were performed in a customized vacuum compatible environmental chamber

Supplementary Note 5b. Test Procedure
A set of strict procedures were followed to ensure consistency throughout the experiments. First, the chamber interior and sample mounting rig were thoroughly cleaned with isopropyl alcohol (IPA) to remove any contaminants. The water reservoir was filled with deionized (DI) water. The heating cables wrapping the environmental chamber were energized to maintain the chamber walls at ≈ 40⁰C. The chamber walls were heated to dry out the inside of the chamber prior to pumping down and prevent condensation during the experiments. The test samples were mounted on aluminum cold plates using PEEK screws and PEEK washers. Then, the test rig with cold plates were connected to the coolant line. Simultaneously, thermocouples were attached to cold plates and chamber walls.
Given the long period of the durability test (> 1 month), a leak test was performed to ensure that no leaks were present before each experimental trial. All diaphragm valves on the degassing chamber and the leak port bellows valve on the environmental chamber were closed. The LN2 trap was filled with LN2. The bellows valve connecting the vacuum pump to the ambient was closed, and another bellows valve connecting the environmental chamber and LN2 trap was opened. Then, the vacuum pump was turned on to initiate the pump down process. The chamber pressure was monitored using the pressure transducer. In the middle of the pump down process, the large capacity chiller was turned on, and the coolant temperature was set to 10°C. The inflow rate of the coolant was monitored using the electromagnetic flowmeter. This process took approximately one hour to achieve the target vacuum conditions (P < 5 Pa). Whenever necessary, the LN2 trap was refilled. The LN2 trap was cleaned midway during the pump down process, since frost would form on the LN2 trap and block the vacuum line during pump down. When the target pressure was achieved, the valve connecting the environmental chamber and the LN2 trap was closed, and the S-53 pump was turned off. Then, the valve connecting the vacuum pump and the ambient was opened to release vacuum in the pump line. The chamber was left under vacuum for >24 hours to perform the leak test, and the chamber pressure was monitored with the pressure transducer. The leak rate was characterized to acquire the fidelity of the data. The initial leak rate was controlled to less than 1 Pa.min -1 , with the long term leak rate controlled to less than 0.1 Pa.min -1 . When the leak test was finished, the vacuum was released by opening the valve connecting the environmental chamber and the ambient.
Once the leak test was finished, the experiment was initiated. The vapor supply diaphragm valve and the water filling valve on the water reservoir were opened, and the tape heater around the water reservoir was turned on with the voltage regulator set to maximum output to boil the water. During the boiling process, the excess water that spilled from the water reservoir was removed. In the meantime, most of the dissolved gas in the water was removed. The temperature of the water and reservoir were monitored using thermocouples. Once the water temperature reaches slightly higher than 100°C for at least 10 minutes, the voltage regulator was turned down, and all the valves were closed. Then, the vacuum pump down process was repeated.
When the target pressure (P < 5 Pa) inside the environmental chamber was reached was collected, and the images of each sample were recorded using the DSLR cameras.

Supplementary Note 5c. Bond Number Calculation to Predict Filmwise to Dropwise Condensation Transition
The droplet Bond number is defined as Bo = f 2 / y 2 , where f is the characteristic lateral length of the liquid droplet, taken to be its final equilibrium radius immediately before departure. The

Durability of HTMS (SAM).
The self-assembled monolayer is obviously thinner than F-DLC. But it is not durable (chemically degrade) because covalent bonding with the substrate becomes chemically unstable when exposed to water. 136,137,138 This is why SAMs have never been demonstrated in the past to be durable dropwise condensation or hydrophobic application coatings.
In our work, we tested samples consisting of Cu and Al tabs coated with SAM of heptadecafluorodecyl-trimethoxy-silane (HTMS), which failed via transition to filmwise condensation after less than one month of testing in pure steam conditions (see Table S4, Supplementary Information). To secure the adhesion between HTMS and substrate, we used the chemical vapor deposition (CVD) method to coat the surface. Specifically, we first ultrasonically cleaned the substrate using acetone for 5 min, followed by rinsing with ethanol, isopropanol (IPA), and deionized water. After drying using a clean N2 stream, we then used air plasma (PDC-001-HP, Harrick Plasma, 30 W, 3-5 min) to clean and activate the substrate. Air plasma cleaning has been shown to increase the number of hydroxy groups on the substrate thus to enhance the HTMS coverage 139 . The plasma-cleaned surface was then put into an atmospheric oven (Thermo Scientific, Lindberg Blue M) together with HTMS-toluene solution (5% v/v) to allow HTMS evaporation and deposition at 80-90ºC for 3 hours.

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The coating methods and protocols are key to secure coating quality. One important aspect is the control of humidity during silane coating. The past studies have highlighted that elimination of water vapor during silane coating can enhance the coating quality. 140,141 In our experiments, we plasma cleaned the sample and reduced the humidity by preheating the beaker and sample in the furnace at ~80ºC for 10-20 min before performing CVD. Although we anticipate the quality of silane coating can be further enhanced by conducting CVD in a vacuum furnace, which has been verified by others, 140,141 our numerous experiments and calculations over the past decade of work demonstrate that these improvements alone cannot overcome the challenges of SAM hydrolysis in steam conditions, which has been well understood and characterized in the past decades.
With our F-DLC multilayer design, we eliminate pinholes, ensuring < 4.2 × 10 -3 (see manuscript for criteria). This design is thermo-mechanically robust and exhibits excellent durability even in humid environments. Consequently, it proves to be a more attractive option compared to SAMs.

Supplementary Note 7a. Droplet Departure Size
We performed condensation experiments focusing on the droplet departure size. Water vapor condensation was performed on both fresh and 3-year durability tested F-DLC surfaces. The duration of the tests on each sample were ~60 minutes. We randomly selected seven droplet departure phenomena for comparison. As shown in Supplementary Figure S8, droplet departure size on durability tested sample is ~28.7% larger than on fresh F-DLC coated surface This is because due to long-term exposure to humid environments, the DI water contact angle of the surface changed from a / r = 99.1°/86.2° to a / r = 70.5°/57.1º.

Supplementary Note 7a. Condensation Heat Transfer
We conducted additional condensation heat transfer experiments to compare the performance of the 3-year durability tested sample with a fresh FDLC coated sample. These experiments were carried out on flat surfaces, unlike the data presented in Figure 3 of the original manuscript, which used tube surfaces. We chose to compare heat transfer coefficients on flat surfaces because the durability test was conducted on such samples, and we only had access to flat surfaces from the 3- year durability test. We did not conduct any durability testing on tube surfaces. Flat surface condensation heat transfer experiments were conducted in a different vacuum chamber ( Supplementary Figures 9a-b), which consists of a main chamber, a boiler or steam generator and a test section (Supplementary Figure 9b). The test section has a cold plate where the temperature is controlled by circulating cold water through a chiller (Supplementary Figure 9c). Additionally, another copper block in the test section had four thermocouples to monitor the temperature gradient ( Supplementary Figure 9c). We attached the test sample at the end of the copper block and inserted a thermocouple into it to monitor the surface temperature.
A set of strict procedures were followed to maintain consistency across experiments and to ensure repeatable results. The test sample was connected to the copper block with thermal interface material and screws, and all connections were checked before the chamber was sealed to ensure there were no leaks. Once the environmental chamber front flange door was sealed, the vapor generator was completely filled with approximately ~5 liters of DI water and heated. After the DI water boiled (>98°C) for at least 20 minutes, the vacuum-pump-down procedure of the environmental chamber was initiated. During the vacuum-pump-down, the pressure inside the chamber was monitored using both installed pressure transducers. This process took ~45 minutes to achieve the target vacuum conditions. The environmental chamber was evacuated to a pressure S-59 of < 5 ± 0.025 Pa with a leak rate of 0.1 Pa.min -1 after chamber isolation. Evacuation eliminated NCGs which add diffusional resistance to condensation heat transfer. 126 The final step in the startup procedure was to initiate the chiller which controls the cold plate and sample temperature. A Polyscience chiller (6860T56A270D) was filled and turned on to the maximum recirculating pressure valve setting, giving a flow rate of ≈ 127 ± 2 g.s -1 . The temperature was set at 7 ± 1°C and the coolant was circulated through the flow loop until the test section inlet temperature ( in) ranged between 7 to 8°C, and was maintained for at least 5 minutes to confirm steady state operation. In summary, the following conditions were met before injecting vapor to the environmental chamber: 1) the liquid inside the vapor generator must reach the boiling point and generate vapor for at least 20 minutes, 2) the coolant and test section temperature must be between 7°C and 8°C for at least 5 minutes, and 3) the chamber pressure must be below 1 Pa.
Once these conditions were reached, the valve in the vacuum-line was closed and the vacuum pump turned off. The vent valve on the vapor generator was closed and the valve connecting the vapor generator to the environmental chamber was gradually opened to allow vapor inside the chamber until it reached a desired pressure. All thermocouples, and pressure transducer readings were recorded over a 10-minute period to obtain steady state condensation measurements.
We performed condensation hear transfer experiments on fresh and 3-year durability tested samples and compared their performance. To ensure one-dimensional (1D) heat transfer, we insulated the test section with a custom-built Teflon block. We inserted four thermocouples into the copper cylinder at known distances to provide a one-dimensional temperature gradient (dt/dx).
By multiplying the temperature gradient with the thermal conductivity of the copper, we obtained the heat flux (q"). We calculated the temperature subcooling by subtracting the measured surface temperature from the saturation temperature corresponding to the saturation vapor pressure. We S-60 calculated the condensation heat transfer coefficient by dividing the heat flux with the degree of subcooling. As shown in Supplementary Figure 9d, we observed < 30% reduction in condensation heat transfer coefficient for the 3-year durability tested sample compared to the fresh F-DLC coated copper surface. This decrease in performance is attributed to a slight increase in condensate droplet size resulting from the reduction in surface wettability after 3-year exposure to humid environments. However, we note that the condensation HTC on 3-year tested surface was ~ 74% higher than an uncoated copper sample which underwent filmwise condensation.

Supplementary Note 8a. Thermal Stability
In order to evaluate the thermal stability of FDLC under steady-state temperature conditions, we subjected the samples to a constant temperature in a controlled furnace environment. These experiments were carried out in both air and an inert atmosphere (Supplementary Figure 10). To anticipate the transition from dropwise to filmwise behavior, we determined the Bond number (Bo) by measuring the contact angle at various temperatures following exposure (Supplementary Figure   10).
In real applications, the elevated temperature is not constant, but rather goes through fluctuations of high and low temperatures. Such temperature variations may result in degradation of the coating due to the mismatch of the thermal expansion coefficient between the substrate and the coating layer. To evaluate the coating's performance in such environments, we conducted a thermal cycling experiment using a laboratory test chamber (Lab event, Weisstechnik) (Supplementary Figure   11a). As depicted in Supplementary Figure 11b, for each cycle, the chamber environment is set to increase from 25°C to 175°C, which takes about an hour, followed by an idle period at 175°C for an additional hour. Then, the chamber temperature starts decreasing from 175°C to -25°C, which takes approximately 2 hours, followed by an hour of constant temperature at -25°C, and finally, the temperature increases from -25°C to 25°C, thus completing the first cycle ( Supplementary   Figure 11b). Each cycle requires approximately 6 hours to complete, and we conducted experiments for a total of 120 cycles, which took roughly 30 days. For this study, we chose both copper and silicon wafers coated with F-DLC. For each substrate, we placed six identical samples on a ceramic boat inside the chamber. After 20 cycles, we removed the first sample of each type.
Thereafter, we took out the rest of the samples sequentially at 40, 60, 80, 100, and 120 cycles.

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After each cycling experiment, we measured the apparent advancing ( a ) and receding ( r ) DI water droplet contact angles using a microgoniometer (MCA-3, Kyowa Interface Science). Even after 120 cycles, we did not observe any significant change in the apparent contact angle, indicating no coating degradation (Supplementary Figure 11c). Both silicon and copper coated substrates showed similar performance.

Supplementary Note 8b. Mechanical Abrasion Characterization
The abrasion resistance of the F-DLC coating on the polished silicon substrate was tested using a Taber Reciprocating abrader (Model-5900) (Supplementary Figure 12). A sliding specimen platform moves in a horizontal, reciprocating motion under a stationary abradant (Calibrase® CS-10, Taber Industries). During the test, the abrasion speed was set to 30 cycles per minute with a 1 N pre-load applied to the abradant holding arm. After each desired abrasion test cycles, abradant residuals on the F-DLC sample were removed by sonicating for 5 minutes in ethanol followed by rinsing with DI water. Then apparent advancing contact angle (θa) and receding contact angle (θr) of water on the samples were measured using micro-goniometry (MCA-3, Kyowa Interface Science Co., LTD).

Supplementary Note 8c. Fluorine Distribution Across the F-DLC Coating (Thickness Direction)
To confirm the distribution of fluorine in the multi-layer F-DLC coating, we conducted EDS line scan analysis on a cross section of coating layer. Supplementary Figure 13 shows the fluorine signal intensity along the scan line. In the Ti, DLN and DLC layers, the signal is nearly zero.
However, the signal intensity in the f-DLC layer increases from bottom of that individual layer to the surface top. S-65

Supplementary Note 9. Effect of Vapor Pressure and NCGs on Coating Condensation Lifespan
Non-condensables (NCGs) are gases that will not condense into a liquid within regular operating conditions. Even a small fraction of NCGs (below 0.1%) seriously hampers steam penetration to the target surface. 142 In condensation applications, there exists a complex vapor concentration profile due to the presence of NCGs. 143,144 The S-67

Supplementary Note 10. F-DLC Thickness and Thermal Resistances
The rational design of the multi-layer F-DLC coating was guided by our physics-based understanding of condensation-induced blistering. The quantitative parameter that describes blistering, , demonstrates that delamination of a hydrophobic coating will occur if > 1. 145 Specifically, the blistering parameter is governed by the pinhole size, d , the base radius of the pinhole-adjunct delaminated region, 0 , the liquid-vapor surface tension of the working fluid DLN. The added DLN layer is well-adapted in conventional DLC multilayers to: 1) enhance adhesion with the Ti layer by silica infusion, 2) provide good thermal stability and act as a stress reliever, and 3) further decrease the pinhole density.
As we discussed above and, in the manuscript, our multi-layer architecture has several functionalities such as: strong interfacial toughness, thermal buffering, mechanical robustness, low surface energy, and pinhole prevention. The current design and thickness are optimized considering all these factors. Reducing the thickness or layer numbers with the existing recipe is not possible, because it will force to sacrifice the durability/robustness.
Furthermore, if we calculate the thermal resistance of the coating at the current design thickness, and then compare it to the thinner coating resistance, we see that both values are substantially smaller than the coolant side thermal resistance, the wall resistance, or the condensation heat transfer coefficient resistance, implying that using a thinner coating is not necessary to enhance performance (Supplementary Table 8). We agree that had the coating been 20 μm thick, the resistance would have been larger and more dominant (Supplementary Table 8).
The overall heat transfer performance from the vapor to the cooling water can be expressed as a series sum of four individual thermal resistances: the internal convective resistance convection , the conduction copper tube wall resistance wall , the conductive resistance imposed by the coating coating , and the condensation heat transfer resistance condensation , as shown in Supplementary

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Considering a copper tube sample having outer diameter OD = 9.53 mm, inner diameter ID =

Payback Period:
To determine the payback period for implementing F-DLC in a power plant condenser, let's consider a typical 500MW power plant with an annual generation capacity of 3.5 TWh.year -1 . 149 Assuming a durable condenser tube leads to a 2% improvement in efficiency, resulting in an overall 2% increase in yearly electricity production, which is equivalent to 0.07 TWh. Based on the previous cost analysis, the total applied coating system cost for F-DLC is ~$4000/m 2 .

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Therefore, the total cost of coating on a 500 MW power plant condenser, which is assumed to have 10 times the surface area of a 50 MW plant, would be approximately (750×10) m 2 ×$4000 ~ $0.3×10 8 .
To determine the payback period, we divide the total cost of coating by the annual benefit from enhanced electrical power generation. Thus, the payback period is approximately ~ $ 0.3×10 8 / $1.15×10 7 /year ~ 2.6 years (excluding the cost of CO2 emissions reduction).
Please note that these calculations consider the provided information and assumptions made.
Actual scenarios may involve additional factors and considerations that could impact the real costs and payback period. It's important to note that the coating cost mentioned earlier is specific to your existing facility. However, when scaling up to a power plant level, it is reasonable to anticipate a significant decrease in costs. Typically, economies of scale come into play, resulting in reduced costs as the size of the operation increases. Therefore, when implementing F-DLC at a power plant level, it is possible that the cost could decrease by a factor of 10 or more compared to the cost at your current facility. This suggests that the payback period for the investment in F-DLC may be even more favorable when considering the cost reduction during scale-up.
Compared to HTMS coating, the applied coating system cost of F-DLC is higher. It is important to consider the long-term benefits of using F-DLC coating, such as its significantly longer service life compared to HTMS coating. The F-DLC system used in this study is performed in a small chamber, in contrast to the widespread use of large and inexpensive PVD systems worldwide.
However, since F-DLC is a combination of PVD and PACVD processes and utilizes the same S-74 facility, scaling up the system for larger production can reduce costs to be comparable to those of PVD systems.
Global high technology markets such as consumer electronics (computer and display) and and used as a protective anti head-disk crash layer. All of these thin films layers are deposited in a single sputtering system with a throughput of 1000 disks per hour and a cost of less than 30 cents per part. 151 Moreover, the equipment used in PVD technique requires low maintenance and the process is environmentally friendly. The coating process does not create any hazardous waste compared to other types of coating, such as electroplating or painting. 152 Moreover, PVD coated products last longer, which reduces the risk of solid waste generation and indirectly helps keep the environment clean. The vacuum environment in the deposition chamber will reduce the gaseous contamination in the deposition process to a very low level. 153 S-75 The global physical vapor deposition (PVD) market size was estimated at USD 23.5 billion in 2021 and it is expected to surpass around USD 47.7 billion by 2030 with a registered compound annual growth rate (CAGR) of 8.18% during the forecast period 2022 to 2030. 2 PVD has been widely adopted in large-scale applications such as solar products, which contribute significantly to the overall market share (Supplementary Figure 15). The increasing adoption of solar energy as a clean energy source propels the demand for solar products. Thus, increasing usage of solar panels and solar cells paves the way for the growth of PVD. 154 The direct method of harnessing solar energy is the solar thermal conversion method using solar absorbers. The absorbers are coated with solar selective coatings often performed with physical vapor deposition (PVD) methods and are used in concentrating solar power (CSP) systems for solar thermal power generation. 155 According to the International Energy Agency (IEA), CSP systems are becoming a crucial technology for mitigating climate change. The IEA report states that by 2050, CSP systems could provide 11.3% of global electricity, with 9.6% from solar power and 1.7% from backup fuels (i.e., fossil fuels and biomass) . 156 At present, parabolic trough technology is the most established and cost-effective large-scale solar power technology for solar thermal power generation. These systems currently have an installed capacity of 870 MW, with 2152 MW under construction and 10 GW in development. As of 2020, more than 4,000 MWe of the worldwide installed operating CSP capacity utilize parabolic trough collectors. 157 The estimated cost of solar collector installation ranges from $90/m 2 -$140/m 2 , 158 which includes manufacturing, purchased components, and installation costs.
CVD is another method of deposition under vacuum and is the process of chemically reacting to a volatile compound from a material to be deposited with other gases, in order to produce a nonvolatile solid that is deposited onto a substrate. This method is sometimes used as pre-coating with S-76 the aim of increasing the durability of the substrates, decreasing the friction, and improving the thermal properties-this means that one can combine deposition methods, like layers of PVD and CVD (Supplementary Table 9), in the same system (similar to our F-DLC deposition method). 113,159 The widespread adoption of PVD technology can significantly reduce costs and carbon footprint, driving demand for PVD equipment in surface coating companies worldwide. 2 Our F-DLC deposition process uses a combination of PVD and PACVD methods in a vacuum system. This implies that the scalability and cost reduction opportunities demonstrated in existing PVD markets, particularly solar collectors, apply to our F-DLC deposition methods. Like PVD, our process is environmentally friendly as it operates within a controlled chamber and generates no solid waste.
However, further research could investigate the use of other chemistries, such as Si, Si+O, or other dopants, to create different types of DLC coatings. This could result in the development of a range of new and improved DLC coatings for various energy applications.
The utilization of scalable deposition methods can be crucial in reducing production costs and enhancing material efficiency, which is essential for sustainable manufacturing processes.
However, it may be questioned whether Dip or Spray coating methods can be used for depositing F-DLC. Unfortunately, our current recipe cannot be applied using Dip or Spray coating. While Dip and Spray coating are generally scalable methods, they present challenges in various ways.
Dip coating is a deposition technique where a substrate is immersed in a liquid solution or suspension of a material and gradually withdrawn at a controlled pace before being thermally treated to cure the coating. This method is scalable, making it ideal for large-scale production.
However, Dip coatings involve solvents that emit volatile organic compounds (VOCs) and perand polyfluoroalkyl substances (PFAS), 160 . PFAS are a class of chemicals that don't naturally S-77 break down, and so they accumulate in water, soil, and in the human body. Studies have shown that high levels increase the risk of cancer and other adverse health effects. 161 The removal of PFASs from surface water, groundwater, soil, sediment and biota is technically extremely difficult and very costly, if at all possible. Considering the effects of PFAs on human health and environment, on February 2023, European Chemical Agency (ECHA) banned around 10000 PFASs. On the other hand, spray coating is another scalable process that can be used on a wide range of materials, from ceramics to metallics. A solution of nanocompounds, adhesives, matrix material, and solvent is atomized by pressure and directed toward the substrate. However, the performance and properties of a coating depend on several parameters, including particle velocity, size, temperature, and position, substrate roughness and chemistry. Spray coatings are directional 162 and difficult to control thickness accurately, 163 and the bind mechanism may not be compatible with complex substrates. Furthermore, achieving a uniform dip or spray coating on a multi-tube condenser poses significant challenges and is almost impossible to accomplish.